Microstructural analysis of wear micromechanisms of WC-6Co cutting tools during high speed dry machining

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Microstructural analysis of wear micromechanisms of

WC-6Co cutting tools during high speed dry machining

Tchadja Kagnaya, Christine Boher, Laurence Lambert, Myriam Lazard,

Thierry Cutard

To cite this version:

Tchadja Kagnaya, Christine Boher, Laurence Lambert, Myriam Lazard, Thierry Cutard.

Microstruc-tural analysis of wear micromechanisms of WC-6Co cutting tools during high speed dry

machin-ing. International Journal of Refractory Metals and Hard Materials, Elsevier, 2014, 42, p.151-162.

�10.1016/j.ijrmhm.2013.08.017�. �hal-00995670�

Microstructural analysis of wear micromechanisms of WC–6Co cutting

tools during high speed dry machining

T. Kagnaya

a,b,

_⁎

_{, C. Boher}

a

_{, L. Lambert}

a,1

_{, M. Lazard}

b,2

_{, T. Cutard}

a

a_{Université de Toulouse, Mines Albi, INSA, UPS, ISAE, ICA (Institut Clément Ader) Campus Jarlard, F 81013 Albi Cedex 09, France}

b_{Université de Lorraine, Laboratoire d'Énergétique et de Mécanique Théorique et Appliquée, LEMTA CNRS-UMR 7563, Ecole Nationale Supérieure des Mines de Nancy (ENSMN),} GIP-InSIC, 27 rue d'Hellieule, 88100 Saint-Dié-des-Vosges, France

a b s t r a c t

Keywords: Turning WC–6Co Tool–chip contact Microstructure Wear mechanism Temperature This
original
study
investigates
the
damages
of
WC–6Co
uncoated
carbide
tools
during
dry
turning
of
AISI
1045
steel
at
mean
and
high
speeds.
The
different
wear
micromechanisms
are
explained
on
the
basis
of
dif-ferent
microstructural
observations
and
analyses
made
by
different
techniques:
(i)
optical
microscopy
(OM)
at
macro-scale,
(ii)
scanning
electron
microscopy
(SEM),
with
back-scattered
electron
imaging
(BSE)
at
micro-scale,
(iii)
energy
dispersive
spectroscopy
(EDS),
X
ray
mapping
with
wavelength
dispersive
spectroscopy
(WDS)
for
the
chemical
analyses
and
(iv)
temperature
evolution
during
machining.
We
noted
that
at
conventional
cutting
speed
Vc
≤
250
m/min,
normal
cutting
tool
wear
types
(adhesion,
abrasion
and
built
up
edge)
are
clearly
observed.
However,
for
cutting
speed
Vc
N
250
m/min
a
severe
wear
is
ob-served
because
the
behavior
of
the
WC–6Co
grade
completely
changes
due
to
a
severe
thermomechanical
loading.
Through
all
SEM
micrographs,
it
is
observed
that
this
severe
wear
consists
of
several
steps
as:
ex-cessive
deformation
of
WC–6Co
bulk
material
and
binder
phase
(Co),
deformation
and
intragranular
microcracking
of
WC,
WC
grain
fragmentation
and
production
of
WC
fragments
in
the
tool/chip
contact.
Thus,
the
WC
fragments
accumulated
at
the
tool/chip
interface
cause
abrasion
phenomena
and
pullout
WC
from
tool
surface.
WC
fragments
contribute
also
to
the
microcutting
and
microploughing
of
chips,
which
lead
to
form
a
transferred
layer
at
the
tool
rake
face.
Finally,
based
on
the
observations
of
the
different
wear
micromechanisms,
a
scenario
of
WC–6Co
damages
is
proposed
through
to
a
phenomenological
model.

1. Introduction

Machining is a largely used process for production of many parts in mechanical industry. Nowadays, the WC–Co cemented carbides are ex-tensively used as cutting tools for their excellent wear resistance, high hardness and good chemical inertness. However, the manufacturing industries use cutting fluids in machining operations which are not en-vironmentally friendly. On the one hand, to avoid the use of these cut-ting fluids, advanced dry machining is performed by manufacturing industries these last decades. On the other hand, in dry cutting, mainly for high cutting speeds, the cutting tool breaks down quickly and dras-tically. This is because high cutting speed leads to high heat generation

and to severe tribological conditions on the tool surfaces, especially on tool rake face.

It is well known that cutting tool wear highly affects the parts' pro-ductivity, production costs and work piece surface integrity. So, a better understanding of the wear mechanisms during cutting remains essential.

Several studies dedicated to machining have been performed in order to understand the cutting tool wear[1–14]. Thus, it was generally found that: wear modes as abrasion, adhesion, plastic deconsolidation, oxida-tion, diffusion and chipping or notching are predominant under various cutting conditions. For example, Komvopoulos[3]reported that abrasive wear is observed on cutting tool when machining AISI 4340. He pointed out that this is due to the combination of high temperature, high work-piece strength, work hardening and chip abrasion. This wear mode has been also reported by other authors[4–7]. Li and Liang[7]found that dur-ing turndur-ing at a cuttdur-ing speed of 50 m/min, AISI 1045 steel adhered to the cutting edge. This adhesion is due to a combination of the ductile work-piece material behavior and of its thermal softening. These results are consistent with Kudou et al.'s [10] one. Furthermore, according to Kudou et al.[10], during high cutting speed machining of AISI 1045, a cra-ter was formed on the tool rake face due to diffusion of cobalt from tool to the workpiece. Diffusion wear is mainly observed at the tool–chip

⁎ Corresponding author at: Université de Lorraine, Laboratoire d'Énergétique et de Mécanique Théorique et Appliquée, LEMTA CNRS-UMR 7563, Ecole Nationale Supérieure des Mines de Nancy (ENSMN), GIP-InSIC, 27 rue d'Hellieule, 88100 Saint-Dié-des-Vosges, France. Fax: +33 3 29 42 18 25.

E-mail address:tchadja.kagnaya@insic.fr(T. Kagnaya).

1_{Now at: CERATIZIT Luxembourg S.a r.l. Route de Holzem L-8201 MAMER,} Luxembourg.

2_{Now at: Institut PPrime UPR 3346, Département Thermique, Fluide et Combustion,} Ecole Nationale Supérieure d'Ingénieurs de Poitiers (ENSIP), Bâtiment B25, Campus Sud, 2 rue Pierre Brousse, 86000 Poitiers, France.

interface and involves chemical reactions between workpiece and tool, because this process is thermally activated. This wear mode is more ob-served at high cutting speeds and essentially for dry machining. Devillez et al.[11]also observed a diffusion wear at rake face of WC–6Co cutting tool when machining AISI 4140 steel under high speed conditions.

This brief review of the cutting tools wear shows that the tempera-ture is one of the most important parameters which influences the tool wear modes. Indeed, M'Saoubi and Chandrasekaran[12] character-ized and identified the various friction zones and wear at the tool–chip interface, based on the temperature measurements by an infrared cam-era. Thus, lots of metal cutting processes are investigated[5,13–18]

(numerically and experimentally) to determine the cutting tool tem-perature. These studies showed that the cutting tool temperature can often reach 1000 °C, especially at tool–chip interface. Consequently, high temperature and severe friction phenomena at tool surfaces deeply modify the mechanical, chemical and thermo-physical properties and can change microstructure of WC–6Co[19–25]. For example Han et al.

[19]show that for temperatures up to 1000 °C, the WC–6Co has a strongly visco-plastic behavior and the accumulation of plastic defor-mation can reach 5%. These observations were confirmed by Östberg et al.[20–22]. According to Childs[23], the oxidation phenomenon of WC–Co takes place at temperatures greater than 500 °C. In addition, in the 900–950 °C temperature range, the mechanisms of dissolution of the tungsten carbide (WC) in the cobalt are activated and change the WC–Co microstructure. Although the mechanics of rock cutting dif-fer to some extent from metal cutting, these observations were recently also confirmed by Beste et al.[24–26]and Angseryd[27]in rock drilling. Thus, Beste et al.[25]showed that the oxidation and the modification of the structure of Co lead to rapid WC–Co wear and a complex mixture of surface damage types, including fragmentation of WC grains.

It is clear that all these phenomena reduce dramatically the wear resis-tance of cutting tool material. Consequently, severe thermal, mechanical and thermo-mechanical solicitations on tool–workpiece and on tool– chip interfaces lead to catastrophic tool wear. Thus, the study of the cut-ting tool wear still remains a challenge in machining processes, since the mechanisms of cutting tool wear in machining, are not yet clearly understood.

Thus the focus of this study is to identify and explain tool wear mech-anisms in relation to the tool material properties and its microstructure. Observations and analyses at macro and micro scales are carried out to clarify complex and combined wear mechanisms of WC–6Co cutting

tool. In addition, microanalyses are performed to determine the chemical composition on tool worn surfaces. Turning tests are carried out using AISI 1045 steel and an uncoated cemented carbide WC–6Co grade. 2. Experiments setup

2.1. Machining experiments

Cutting tests are carried out on a T400 REALMECA CNC turning ma-chine with a spindle speed up to 4000 rpm and a drive motor rated up to 7 kW (Fig. 1).

The workpiece material used throughout all these experiments is an AISI 1045 laminate steel in the form of round bars with 100 mm diam-eter. The mechanical properties, thermo-physical properties and chem-ical composition of workpiece are listed inTable 1. A pre-machining is made previously along the workpiece in order to eliminate any surface defect that could adversely affect the machining results.

A STCGL 2525M16 (40CrMoV12) tool holder with an uncoated insert TCMW 16T304 (WC–6Co) is used. The cutting tool geometry characteris-tics, the chemical composition, and the physical and mechanical properties of this carbide tool grade are summarized respectively inTables 2 and 3. The test conditions are summarized inTable 4. For each experiment, a fresh part of the cutting tool is used and experiments are replicated twice for each cutting conditions in order to minor experimental error. During experimental tests, two thermocouples (K type) with 0.25 mm of diameter, are embedded in the cutting tool in order to mea-sure the temperature evolution during machining process. The first (TC1) and the second (TC2) thermocouples are embedded respectively at 0.5 mm and 1 mm from the rake face of the cutting tool as shown in

Fig. 1(b). The three components of the cutting forces (cutting force Fc, feed force Ff, and depth of cut force Fp) are recorded for all the tests, using a dynamometer (KISTLER table, Model 9257B).

2.2. Characterization methods

Tool wear analyses are performed with two observation scales. Macroanalysis of tool wear is first performed with both optical micros-copy (LOM) and quasi confocal optical scanning microsmicros-copy (COSM) Altisurf[28]to measure the geometrical parameters of the wear surface of the tool. The main advantage of the LOM, is to have a global view of

Fig. 1. (a) Experimental setup for machining tests, (b) detail of cutting tool insert.

Table 1

Mechanical properties and chemical composition of AISI 1045 steel.

Components C Si Mn P S Cr Ni Mo Cu %wt 0.45 0.22 0.66 0.027 0.032 0.26 0.15 0.02 0.18 Yield strength (MPa) 437

Rupture strength (MPa) 704

Table 2

Uncoated insert TCMW 16 T304 (WC–6%Co) geometrical char-acteristics.

Nose radius 0.4 mm

Cutting edge radius 0.04 mm

Rake angle 0°

the cutting tool wear surface, to identify the main wear modes and to measure some parameter of tool wear, as the flank wear width and the tool-chip contact length. However to obtain the crater wear depth and the crater wear volume, COSM Altisurf® 520 apparatus with an op-tical probe having a precision of about 2 nm/mm is carried out. The 3D rake face surface data processing is performed with the Altimap analysis software which offers a wide range of analysis and measures.

The observations of tool wear at the micro-scale are investigated with a scanning electron microscope (SEM) with back-scattered electron im-aging (BSE). Energy dispersive spectroscopy (EDS) and X-ray mapping analysis with wavelength dispersive spectroscopy (WDS) are performed in order to identify the chemical elements at the surface of the cutting tool. All EDS spectra are performed using 20 kV accelerating voltage and X-ray mapping with WDS is complementary to EDS because it gives a bet-ter resolution for light elements as oxygen. All these analyses are useful for a better understanding of damage WC–6Co micromechanisms. 3. Results and discussion

3.1. Cutting forces and temperature analysis

Fig. 2(a) presents the evolution of the cutting forces, generated dur-ing the AISI 1045 machindur-ing, versus the cuttdur-ing speed. Fc, Ff and Fp are respectively the main cutting force, the feed force and the depth of cut force. The force component (Fig. 2) shows a decreasing trend with an in-crease in cutting speed, especially beyond 250 m/min. This evolution of cutting forces is in good agreement with general understanding in metal cutting[8]. One can notice that this reduction is particularly important with the feed force. Indeed,Fig. 2(b) shows that the specific cutting en-ergy, SCE, calculated by Eq.(1)varies slightly according to the cutting speed, while the tool–chip friction heat flux, Qf, obtained by Eq.(2)

increases as cutting speed increases.

SEC¼tc_VVc Fc

mat ð1Þ

Qf¼ Vchip F2fþ F2p

! "1=2

ð2Þ where Vc is the cutting speed (m/s), Vmatis the removed chip volume

(m3_{), t}

cis the cutting time (s), and Vchipis the chip flow velocity (m/s).

Therefore, the cutting force reduction can mainly be attributed to the softening of the AISI 1045. A higher cutting speed involves a higher heat flux and so a higher temperature (Figs. 2(b) and3). Thus, a high heat flux generated at high cutting speeds produces morphology changes of the chip and so, the outstanding feed force reduction is due to the chip segmentation[14].

To take into account the temperature in tool wear analysis, the cut-ting tool temperature is measured through two isolated K-type thermo-couples (ϕ = 0.25 mm).Fig. 3(a) and (b) shows the evolution of the cutting tool temperature (thermocouples TC1 and TC2), versus to cut-ting time and for four cutcut-ting speeds.Fig. 3depicts that the temperature increases with increasing cutting speed and the reached temperature (about 600 °C) is enough high to influence cutting tool wear.

For example, the highest temperatures recorded by the thermocou-ple nearest to the rake face (TC1), for the cutting speeds 100 m/min and 400 m/min after about 30 s of machining time, reach respectively of 400 °C and 820 °C. These temperatures are high enough to modify the tool material behavior and the microstructure of WC–6Co.

According toFig. 3, it is observed that when the cutting speed “Vc” is lower than 300 m/min, the TC1 and TC2 temperatures become constant respectively 5 s and 10 s after the start of the experiment. On the other hand, for Vc = 400 m/min, TC1 and TC2 temperatures are not stabi-lized after 25 s of machining. Based on the TC1 and TC2 positions, the differences of the temperature evolutions between Vc≤ 300 m/min and Vc = 400 m/min are due to the various heat input in the tool– chip interaction areas (Fig. 2(b)). This differences can also be due to the increase of the crater wear for Vc = 400 m/min, i.e. a fast decrease of the distances between the thermocouples and the tool rake face and consequently the rapid increase of the recorded temperatures. These temperature measurements explain that two wear modes exist: a pro-gressive wear mode for cutting speeds Vc≤ 300 m/min and a fast wear mode for cutting speeds Vc N 300 m/min. This conclusion is con-fi_{rmed inSection 3.2.1“Macroscopic scale tool wear analysis”.}

As mentioned above, temperature considerably impacts the cutting tool wear. However, direct temperature measurements at the tool– chip and tool–workpiece interfaces are very difficult to get and even impossible because these interfaces are unattainable during machining

Fig. 2. (a) Cutting forces evolution, (b) specific cutting energy and tool–chip friction heat flux vs. cutting speed (f = 0.1 mm/rev, doc = 1.1 mm).

Table 3

Thermo-physical and mechanical properties of the cutting insert (WC–6Co). Thermal conductivity Temperature (°C) 20 100 200 300 500 600

Value (W/m °C) 117 110 97 86 85 83 Specific heat (J/kg °C) 222

Mean coefficient of thermal expansion (μm/m °C) 6.30 Density, ρ (kg/m3_{) 14900} Young modulus (GPa) 600 Transverse rupture

strength (MPa)

1620 Hardness Hv10(MPa) 1550

Table 4

Cutting test conditions.

Cutting speed 100≤ Vc (m/min) ≤ 500 Constant feed rate f = 0.1 mm/rev Constant depth of cut doc = 1.1 mm

process. So, the temperature distribution at tool–chip interface, obtained by numerical simulation and discussed in previous studies

[14,29], is used to well analyze progressive tool-wear. Some results of these previous studies are presented here in order to better establish the correlation between temperature levels and WC–6Co damages.

Fig. 4(a) presents the definition of the tool–chip contact area on which estimated values of friction heat flux, generated at chip–tool inter-face for each cutting condition, are applied to solve heat transfer numeri-cal simulation.Fig. 4(b) and (c) illustrates an example of the excessive temperature distribution with the high temperature gradient on tool–

chip contact for the following cutting conditions: Vc = 400 m/min, f = 0.1 mm/rev and doc = 1.1 mm. For this example, the ultimate inter-face temperature varies between 1280 °C and 650 °C. As it can be seen in

Fig. 4(b), two temperature profiles are plotted from the cutting tool edge to the end of tool–chip contact. The first profile corresponds to line 1, which is located at 0.3 mm from the tool tip and the second one corre-sponds to line 2, which is located at 0.9 mm from the tool tip. According toFig. 4(c), the temperature along line 1 decreases from 1150 °C at the tool edge to 900 °C, while the line 2 temperature decreases from 980 °C down to 760 °C. Using these numerical simulation results, the average

Fig. 4. Cutting tool: (a) definition of tool–chip contact area, (b) rake face temperature distribution, (c) temperature profiles on the rake face[14,24](Vc = 400 m/min, f = 0.1 mm/rev, ap = 1.1 mm, after end of cut).

temperatures on tool rake face are calculated and are respectively 891 °C, 1002 °C and 1082 °C for cutting speeds of 250 m/min, 300 m/min and 400 m/min.

3.2. Cutting tool wear analysis

Fig. 5illustrates progressive wear of the crater (rake face) and the fl_{ank face according various cutting times and cutting speeds (100–} 250–300–350 and 400 m/min). These observations are in good agree-ment with time–temperature curve evolutions observed previously fromFig. 3. For each condition, tool–chip contact length, Lc (crater wear width) and flank wear width VB are measured. Lc and VB values increase according to cutting times for all cutting speeds.Fig. 5also sug-gests that for the same time (for example time = 10s), increasing cutting speed induces an increase of VB. Instead, Lc is not a pertinent pa-rameter since tool–chip contact length depends strongly on the compe-tition between the chip curvature, the chip flow velocity and the local softening of the chip material.

At high speed machining, the wear of the tool rake face predomi-nates and therefore tool life is determined by the crater wear depth

until the edge failure occurs. For a good understanding of tool wear evo-lutions, mainly in terms of crater wear, a typical crater wear depth “KT” and crater wear volume “VKT” are measured with COSM.Fig. 6 (a) shows an example of 3D crater wear topography measured by COSM for Vc = 500 m/min.Fig. 6 (b) illustrates the crater wear contour used to calculate the crater wear volume “VKT”.Fig. 6(c) illustrates one crater wear profile along the x axis which is perpendicular to the cutting edge according toFig. 6(a) and so, the values of crater wear depth “KT” in plane (x, z) can be measured.Fig. 7shows the evolution of the KT and VKT values versus cutting speeds. For Vc≤ 250 m/min, both KT and VKT are close to zero, however, above Vc = 250 m/min, the increase of cutting speed induces the increase of both KT and VKT.

As previously mentioned, crater wear is more important at high cut-ting speed machining and only worn surfaces of the rake face were ex-amined by integrating the temperature distribution at the tool–chip contact (Fig. 4). To extend our study, investigations from SEM micro-graphs are carried out. Therefore, macroscale and microscale tribologi-cal behaviors between tool and chip are studied from three wear zones as showsFig. 8(a) and (b) through SEM observations of the surfaces and cross-sectionsFig. 8(c). The analysis of WC–6Co cutting

tool wear and discussion of each zone are presented separately includ-ing SEM observations and EDS analyses to evaluate adhesion of work material on worn surfaces of the cutting tool.

3.2.1. Wear analysis in zone 1

This small zoneFig. 8(a), near to the cutting edge, corresponds to the sticking zone. In that zone, the contact surface appears smooth

and clean, which induces quasi-static friction, (sliding velocities are close to zero). This interpretation is confirmed by Zorev results[25]

that identified this zone like an intimate contact tool–chip zone. The size of this zone is strongly affected by cutting speed and determines the position of the maximum crater wear depth, KT. The increase of the cutting speed leads to the decrease of the sticking zone. The size re-duction of the sticking zone impairs the tool edge resistance. Then for high cutting speed, crater wear controls the cutting edge life. Higher normal and shear stresses (Fig. 8(a)) in this zone are involved in the severe damage mechanisms.

Several damage modes are observed. The microstructure of WC–6Co appears disorganized (Fig. 9(b), compared to the initial oneFig. 9(a)). This disarrangement is the result of local intergranular microcracks at WC/WC or WC/Co interfaces or chipping due to the alternate stress and to the impact stress leading to particle release. One can note that microcracks extend at a distance equivalent to several grain sizes as showsFig. 9(c). In addition, plastic deformation steps of WC grains resulting from intragranular slips and probably according to crystallo-graphic plans of WC, are observed on crater worm surface inFig. 9(d) (shown by slip lines on the surface of WC grains). Consequently, when the plastic deformation of WC is important, the intragranular microcracks could appear. But intragranular microcracks can also appear when a stress concentration exists at WC/WC. This wear mech-anism was also observed by Beste et al.[25]in rock drilling. An example of intragranular microcrack is illustrated inFig. 9(e). The void observed inFig. 9(f) shows that the grains and/or the fragments of WC are time to time torn off from tool–chip contact. A small transferred layer is also ob-served in zone 1 and it represents WC–6Co/(chip material) weld junction.

Fig. 6. (a) 3D crater wear topography (Vc = 500 m/min, feed = 0,1 mm/rev et doc = 1,1 mm), (b) determination of crater wear volume (VKT, μm3_{), (c) profile of crater perpendicular to} cutting edge, determination of KT (μm) in plane (x, z).

3.2.2. Wear analysis in zone 2

When the chip leaves the zone 1, its velocity increases progressively up to a steady chip flow velocity. Therefore zone 2 corresponds to an in-tense tool–chip friction zone, in which the chip flow velocity is stable. The shear stress is optimum in this zone (Fig. 8(a)). However due to the severe tribological conditions and to the high temperature levels at the tool–chip interface in this zone, the position of the maximum cra-ter wear depth, KT is located in zone 2.

Damage micromechanisms observed in zone 2 are the same with those described for zone 1 except the process of disarrangement of the WC microstructure. However the phenomena of intergranular and intragranular microcracking as well as wrenching of WC grains are less important. The observation ofFig. 10(a) shows several abrasive micro-scratches or slip line traces on the surface of WC grains in the chip flow direction. This second zone exhibits an important adhesion layer (Fig. 10(b)) due to a combination of the high friction forces be-tween the tool and the chip and the high temperature level in this zone. After its formation, adhesion layer could be periodically evacuated by the chip. This wear phenomenon can be summarized like following periodic phenomena of “formation–stacking (growth of adhesion layer)–pulling out”. During adhesion layer elimination, more or less WC grains can be debonded and the presence of the WC grain in adhe-sion layer (Fig. 10(b)) illustrates this remark. These wear debris, i.e. WC grain fragments, could increase the abrasion phenomena on rake face, during the machining process.

3.2.3. Wear analysis in zone 3

It is a crown shape zone, which is surrounded by zone 2 (Fig. 8(a)). It characterizes the entry and the exit of the intense friction zone (artificial secondary rake face and artificial backwall contact zones). In this zone, adhesion of the machined material is more important. When the chip leaves the zone 1, the shear stress starts to decrease and reach an optimum. Towards the exit of the chip, under the effect of the contact pressure and chip curvature, an unsticking resistance of the chip

appears which involves an increase in the shear stresses before these stresses relapse to zero when the chip loses its contact with the tool.

In zone 3, an important adhesion layer is formed as shown inFig. 11. The curvature and intense shear deformation of the chip in this zone in-duce chip ploughing phenomenon and so, can explain this important adhesive layer. The accumulation of workpiece material in this zone modified strongly the parameters of the cutting rake face geometry (rake angle for example).

3.2.4. Complementary EDS and X-ray mapping analysis

According to some studies[10,30–32], diffusion wear mechanism can occur at tool–chip interface when machining free cutting steels be-cause Co and W elements can move towards chip material and Fe element can diffuse into WC–6Co material.

In order to investigate the effect of diffusion wear mechanism on the crater wear process, EDS analysis is performed in small zone (ZA) as it can be seen in the cross-section ofFig. 11(c) and the result of the EDS analysis is shown inFig. 12. Except for oxygen element, all the chemical components detected in the adhesion layer are the same ones as those of the workpiece material. The high oxygen content indicates that the adhesion layer is oxidized, as a consequence of the elevated tempera-ture generated at tool–chip according toFig. 4(b). EDS results indicated that Co and W elements have not diffused into the adhesion layer.

In certain circumstances, EDS results may contain peak overlaps for iron and cobalt components. So 2D element maps with WDS which can detect trace elements are very useful for displaying element distri-butions in sample.Fig. 13shows the spatial distribution of four elements (Fe inFig. 13(a), O inFig. 13(b), W inFig. 13(c) and Co inFig. 13(d)) according in square zone defined by dotted lines onFig. 11(c). From qualitative point of view, the high concentration of white pixels in the images indicates a high content of the considered element. The position of (adhesion layer)/(WC–6Co) interface is materialized by a dotted lines. In the adhesion layer, the higher content of Fe compared to O, W or Co can be observed, and the presence of oxygen confirms that the

Fig. 8. (a) Evolution of normal and shear stress at tool–chip contact zones (feed = 0.1 mm/rev, doc = 1.1 mm), (b) three zones define for wear analysis and (c) SEM micrographs of crater wear of the cutting tool cross-section, perpendicular to the cutting edge (Vc = 400 m/min, feed = 0.1 mm/rev, doc = 1.1 mm).

Fig. 10. Crater wear SEM micrographs for detail observations: (a) zone 2 inFig. 8b, (b) zone A and (c) zone B inFig. 8c. Fig. 9. (a) Initial micrograph of WC–6Co, (b)–(f) crater wear SEM micrographs for detail observations zone 1inFig. 8b.

adhesion layer is oxidized. As observed in element maps shown in

Fig. 13(c), bottom part of dotted line zone contains a large amount of tungsten and the concentration of Co can be considered as insignificant compared to W (Fig. 13(d)).

Through chemical analyses, it is not obvious to conclude that there is diffusion of the tool elements into the adhesion layer (the chip) and vice-versa. However, the diffusion wear phenomena takes perhaps place in the thin adhesion layer at tool–chip interface and consequently requires other more powerful techniques than those carried out in this study.

A qualitative summary of the WC–6Co damage mechanisms, which were identified through SEM, EDS and WDS analyses is presented in

Table 5. ThereforeSection 3.2.5presents complementary phenomeno-logical models suggested to detail the micromechanisms that control the wear behavior mentioned above.

3.2.5. Discussion of the phenomenological model

It is important to note that during the machining process, numerous complex and intensive interactions occur in the cutting zone, mainly at the tool–chip interface. These interactions result in different phenomena like friction, plastic and elastic deformations, chemical reactions and heat generation. These solicitations become more severe in dry high speed machining and lead to a faster tool wear. On the basis of observa-tions carried out in this present work, the aim of this section is to pro-pose model to explain the damage mechanisms of the WC–6Co.Fig. 14

presents this phenomenological model that has been established from the tool wear mechanisms identified mainly in zone 2 of crater wear. 3.2.5.1. Stage 1.Under severe friction conditions (high shear stresses) at the tool–chip interface, the local temperature rises up to several hun-dreds of degrees (about 1280 °C—Fig. 4(b)). Referring to the results of Östberg et al.[20], such a temperature level can lead to the plastic deformation of WC–6Co, because plastic deformation transition of the WC–6Co appears towards 900 °C. Moreover at such temperature level, the binder phase, which is a Co–W–C alloy, becomes very ductile since its melting point is lower than that of pure cobalt (1500 °C). Thus, the binder phase between the WC grains can be extruded or partially re-moved, due to the movement of the WC grains at the crater surface. These phenomena involve a rearrangement between the WC grains and the binder phase as that is observed inFig. 9(b) and can start the debonding mechanism of the WC grains as schematized byFig. 14(a). 3.2.5.2. Stage 2.Locally, when the binder phase between the WC grains is removed, and hard-on-hard WC/WC contacts are established, an in-crease of the stress level in the WC grains occurs. Taking into account the angular shape of the WC grains, stress concentrations appear at angular contact points as shown inFig. 14(b).

3.2.5.3. Stage 3.Due to stress concentrations, temperature and cyclic me-chanical loadings during machining, the intragranular deformation mechanisms of WC grains are activated as observed inFig. 8(c) and as shown schematically inFig. 14(c). The plastic deformations observed

Fig. 11. Crater wear SEM micrographs for detail observations: (a) and (b) zone 3 inFig. 8b, (c) zone C inFig. 8c.

in WC grains are in good agreement with the results of literature

[19,20,33]. Furthermore, such intragranular phenomena can initiate microcracks. Furthermore, the decrease of WC hardness above of 800 °C[34], can lead to intragranular propagation of the microcrack after its initiation. This mechanism produces fragments of WC grains as observed previously in the underlying layer at the friction surface in

Fig. 10(b).

3.2.5.4. Stage 4.When the WC fragments are produced, they are embed-ded in chip or in the adhesion layer and they take part both in abrasive wear mechanism at tool rake face and in the debonding process of the WC grain, as shown inFig. 14(e)–(f). Moreover, towards the tool– chip separation zone, especially strong adhesion layer is formed. It is due to the fact that in this zone both mechanical loads (normal pressure and shear stress) and the temperature levels are less important com-pared to those of zones 1 and 2.

4. Conclusions

This work provides a better understanding of the different damage at the micro-scale of WC–6Co in machining processes. The main conclu-sions of the present work can be summarized as follows:

• The cutting tool temperature increases with the cutting speed, while cutting forces decrease as cutting speeds increase. Increasing the cut-ting tool temperature leads to decrease the friction coefficient and consequently the cutting pressure.

• The cutting tool wear depends on cutting speed. At conventional cutting speeds, a normal wear of the flank tool is observed. For high cutting

speeds, a faster wear rate on the rake face is predominant. At a macro-scopic scale, adhesion, abrasion and chipping wear are observed. The crater wear mode is dominant during machining AISI 1045 at high cut-ting speeds with WC–6Co cemented carbide cutcut-ting tools.

• The catastrophic wear mechanism of WC–6Co tools during high-speed machining of AISI 1045 is activated by the coexistence of two main factors: severe tribological conditions on cutting tool and heat generation. These conditions lead to the following tool wear micromechanisms:

- Disarrangement of the WC–6Co macrostructure due to the severe tri-bological conditions at tool–chip and tool–workpiece interfaces. This disarrangement leads to microcracks at WC/WC or at WC/Co inter-faces or chipping due to the alternative stress and to the impact stress.

- Plastic deformation and intragranular microcraking of the WC grains. This intragranular microcraking leads to the fragmentation of WC grains and to the production of WC debris at the tool–chip interface. - According to SEM observations, there is adhesion at the tool–chip separation zone and this adhesive layer is formed according to a “for-mation–stacking–pulling out” process. During machining, it is re-moved periodically by the chip and this removal can be amplified by the debonding of WC fragments.

• From the chemical analyses carried out in the zone of strong adhesion, it is not obvious to conclude that there is diffusion of the tool elements in the adhesion layer (the chip) and vice-versa. However, the diffusion wear phenomenon perhaps takes place in the thin adhesion layer at tool–chip interface and consequently requires other techniques than those used in our investigations.

• Finally, a phenomenological model is proposed in order to identify the different wear sequences of the WC–6Co tool during dry turning process.

Acknowledgments

Authors thank ACTARUS S.A.S. Company and European Research Center for Rapid Prototyping and Rapid Tooling (CIRTES) for supporting this work.

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